Investment shell cracking

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Investment shell cracking Investment shell cracking

Edward A. Druschitz

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i

INVESTMENT SHELL CRACKING

by

EDWARD ALAN DRUSCHITZ

A THESIS

Presented to the Faculty of the Graduate School of the

MISSOURI UNIVERSITY OF SCIENCE AND TECHNOLOGY

In Partial Fulfillment of the Requirements for the Degree

MASTER OF SCIENCE IN MATERIALS ENGINEERING

2009

Approved by:

Dr. Von L. Richards, Advisor Dr. Ronald Kohser

Dr. Frank Liou

ii

iii

PUBLICATION THESIS OPTION

This dissertation consists of three articles submitted for publication as follows. Each

article is prepared according to the respective style of the publication. Pages 3-17 have

been submitted for publication in the International Journal of Cast Materials. Pages 18-

32 have been published by The Minerals, Metals, and Materials Society. Pages 33-45

have been published by the Investment Cast Institute.

iv

ABSTRACT

Shell cracking is the single greatest problem affecting investment casters. A

clearer understanding of the factors affecting the melt profile of the wax can be gained

using computational fluid dynamics (CFD) to model the interaction among 1) the thermal

conductivity of the wax, 2) the thermal conductivity of the shell, and 3) the temperature

of the autoclave during the autoclave de-waxing cycle. The most favorable melt profile

results from a high autoclave temperature (438°K to 458°K) and saturated thermal

conductivity of the shell (1.36 to 1.40 Wm-1k-1) in conjunction with a low wax thermal

conductivity (0.33 Wm-1k-1). These parameters reduce the likelihood of shell cracking as

a result of wax bulk expansion.

Thin wall ice patterns can be invested by coating patterns with a tridecane

interface agent, using a 20 wt% ethyl silicate binder with 20 wt% fiber-containing fused

silica flour added after the primary coat, delaying the application of the catalyst by 4

hours, and maintaining a -10°C environment. A -10°C or lower environment increased

shell strength and improved surface finish. It also found that tridecane resulted in a 7%

loss of thickness in ice 3.175mm (0.125 inches) and thicker.

Oxidation along the leading edge of cast Fe-15Cr-4.5Ni-3Cu (15-5 PH) stainless

steel marine propellers leads to costly non-value added finishing. The addition of an

extra seal coat of slurry after autoclaving and a slower cooling rate can reduce oxidation

by 1.7 to 1.4mm2mm-1 of oxidation per millimeter of blade length.

v

ACKNOWLEDGMENTS

First and foremost, I would like to thank Dr. Von Richards; without him none of

this would have been possible. He has been more patient then I could ever have hoped as

I have stumbled to gain a foothold in metallurgy. His time and patience have been

greatly appreciated during my struggle to understand materials engineering.

Next, I would like to thank my committee members, Professor Ronald Kohser and

Professor Frank Liou for their guidance and time. They have helped me gain a

fundamental understanding of key issues important to this research.

I would also like to thank my family, Alan, Lori, and Laurel. Without my father’s

constant guidance, patience, and undying support I may not have seen this through to the

end; he is a constant inspiration for me. He is my role model as a developing professional

in the metallurgical field. My mother and sister have been constant pillars of support,

and I cannot thank them enough.

In addition, I would like to thank a few professors who have also helped me

greatly during my time here: David Van Aken, Greg Hilmas, Jeffrey Smith, William

Fahrenholtz, and Mary Reidmeyer, your constant words of encouragement have helped

me more then any of you know.

I would also like to thank a few graduate students who have helped me stay on

track and at times reminded me to never give up: Hank Rawlins, Ryan Howell, and

William Dewy Peach, you have truly helped me find myself as a graduate student and a

person.

vi

TABLE OF CONTENTS

Page

PUBLICATION THESIS OPTION .............................................................................. iii

ABSTRACT…………………………………………………………………………...iv

ACKNOWLEDGMENTS .............................................................................................. v

LIST OF ILLUSTRATIONS ...................................................................................... viii

SECTION

1. INTRODUCTION ........................................................................................ 1

PAPER

I. Parametric Modeling of the Autoclave De-waxing Process ....................................... 3

1. ABSTRACT .................................................................................................. 4

2. INTRODUCTION ........................................................................................ 4

3. BOUNDARY CONDITIONS AND PARAMETERS ................................. 8

4. COMPUTATIONAL PROCEDURES ....................................................... 12

5. RESULTS ................................................................................................... 14

6. DISCUSSION ............................................................................................ 15

7. CONCLUSIONS......................................................................................... 16

8. FUTURE WORK ........................................................................................ 16

9. REFERENCES ........................................................................................... 17

II. Investment Shell Building on Ice Patterns ............................................................... 18

1. ABSTRACT ................................................................................................ 19

2. INTRODUCTION ...................................................................................... 19

3. EXPERIMENTAL PROCEDURE ............................................................. 24

3.1. PATTERN LOSS ................................................................................ 24

3.2. DELAYED CATALYST APPLICATION ......................................... 25

4. RESULTS ................................................................................................... 26

4.1. PATTERN LOSS ................................................................................ 26


5. DISCUSSION ............................................................................................. 30

5.1. PATTERN LOSS ................................................................................ 30

vii


6. CONCLUSION ........................................................................................... 30

7. FUTURE WORK ........................................................................................ 31

8. REFERENCES ........................................................................................... 32

III. Oxidation During Solidification of 15.5 PH Marine Propellers ............................. 33

1. ABSTRACT ................................................................................................ 34

2. INTRODUCTION ...................................................................................... 34

3. EXPERIMENTAL PROCEDURE ............................................................. 40

3.1. SHELL STRENGTHENING .............................................................. 40

3.2. COOLING RATE EFFECTS .............................................................. 42

4. RESULTS ................................................................................................... 42



5. DISCUSSION ............................................................................................. 43



6. CONCLUSION ........................................................................................... 44

7. FUTURE WORK ........................................................................................ 44

8. REFERENCES ........................................................................................... 45

SECTION

APPENDIX…………………………………………………………………………...46

VITA………………………………………………………………………………….68

viii

LIST OF ILLUSTRATIONS

Figure Page

PAPER I 2.1. Schematic of heat transfer in the autoclave. Steam condenses on the shell’s surface and saturates the shell. Heat is transferred through the shell to the under lying wax, resulting in melting [Gebelin, 2001]. ......................................................... 6 3.1. Wax block surrounded by ceramic investment casting shell with boundary condition Tw (temperature at wall). .............................................................................. 9 3.2. Melt profiles of the wax as a result of the calculated distribution of the liquid fraction at 480s (left) and 640s (right). ...................................................................... 12 PAPER II 2.1. Rapid freeze prototyping (RFP) machine inside a deep freezer. Water is deposited on the liquid nitrogen chilled substrate, which is moved by the XY table. 20 2.2. Schematic of slurry tank system. It is driven by an electric motor with built-in

reducing gears (16rpm) rotating a five gallon bucket using a ½ inch drive belt. ....... 22 2.3. Step plate test article used for ice pattern, thin wall, investment casting trials (units are inches). ....................................................................................................... 23 2.4. Diagram of shell coats: The primary coat determines surface finish, detail coats build pattern detail, backup coats provide strength, and the seal coat holds the last backup layer’s stucco in place. ...................................................................... 23 4.1. Ice pattern loss and resulting shell cavity thicknesses. Samples one through six were tridecane coated, samples seven, ten, and eleven were not coated. Delamination of the primary coat caused cavities to show a net increase in thickness.

.................................................................................................................................... 27 4.2. Percentage loss based on step thickness. Samples one through six (tridecane coated) experienced minimal loss of thickness compared to non-coated step plates. .................................................................................................................. 28 4.3. Investment shell exhibiting primary coat delamination. ............................................ 28 4.4. Step plate shells with and without tridecane coating. Tridecane coating increased retention of the primary coat. Step plate thickness increases from left to right. ....................................................................................................................... 29 PAPER III 2.1. Oxidation along the leading edge of a cast marine propeller. .................................... 34

ix

2.2 Effect of particle shape and volume fraction on fracture toughness [Richardson,

2006]. ......................................................................................................................... 36 2.3. Diagram of shell coats, the primary coat determines surface finish, detail coats build pattern detail, backup coats provide strength, and the seal coat holds the last backup layer’s stucco in place. ...................................................................... 37 2.4. Illustration of cooling rate based on position. The inside propeller blades are

surrounded by hot castings and will therefore cool slower. ....................................... 38 2.5. Thermal image of solidifying casting; areas in red indicate a temperature of 428°C (802°F). ....................................................................................................... 39 2.6. EDS image of leading edge oxidation of cast 15-5PH Stainless. .............................. 40 3.1. Oxidation on leading edge of propeller blade. ........................................................... 41 3.2. Image analysis representation of oxidation on leading edge of propeller blade. ....... 41

x

LIST OF TABLES Table Page

PAPER I 3.1. Parameters used to determine the influence of shell and wax conductivity and

autoclave temperature on the melt profiles of wax during autoclaving. .................... 10 3.2. Additional properties of the wax and shell required for the completion of the

parametric simulation. ................................................................................................ 11 5.1. Times required for melting given varying shell and wax conductivities and autoclave temperatures. Increased autoclave temperature, saturated shell

conductivity, and low wax thermal conductivity resulted in the highest DT time of 48 seconds. ..................................................................................................... 14 PAPER II 4.1. Four-point bend test results for delayed catalyst application. A four hour delay

resulted in the highest average strength (16.4 N/cm2). .............................................. 29 PAPER III 4.1. Oxidation per unit length of blade (mm2mm-1). ......................................................... 42 4.2. Oxidation (mm2mm-1) as a function of position ........................................................ 43

1. INTRODUCTION

Investment casting can be considered either a new or an old technology,

depending on one’s perspective on industrial history and genealogy. Industrial

investment casting began with the need for intricate turbine blades during World War II.

However, the process of lost wax dates back to pre-Christian Egypt and Chinese

dynasties as early as 4,000 BC. Older pots, vases, wine goblets, and religious artifacts

display intricacy achieved using lost wax. Older methods of investing the wax involved

packing clay around bee’s wax patterns before firing, creating castings devoid of parting

lines. However, advances in ceramics and shell building dramatically changed the

investment casting industry. Corning Glass Works patented and marketed a technique

called Glascast in 1957; simultaneously, Watertown Arsenal introduced a process called

sintered alumina mold. Both processes are recognized as the original ceramic shell,

nonflask, investment casting technique [AFS, 1993]. Investment casting begins with the

creation of a wax part by injection. The part is assembled into a tree wherein numerous

parts share a single downsprue. The tree is invested and stuccoed before removing the

wax using an autoclave or boilerclave. The shell is fired to add strength before filling

with molten metal. The metal is allowed to cool before removing the castings for

finishing.

The primary goal of this research was to reduce or eliminate shell cracking in

investment castings. This project involved: 1) continuing efforts to develop a predictive

parametric model of autoclave dewaxing, since cracking often occurs in the autoclave

process, 2) building thin wall shells for the casting of aluminum metal matrix composites

using ice patterns in association with rapid freeze prototyping (RFP) technology, and 3)

2

reducing cracking during solidification after pouring, resulting in increased casting

quality in cast 15-5 PH marine propellers.

Parametric studies on heat flow and melt front progression should be conducted to

determine real world boundary conditions via computer modeling of autoclave

temperature, shell thermal conductivity, and wax conductivity.

Many problems arise from the volumetric expansion of wax during the autoclave

de-waxing cycle. Extensive research has sought to alleviate or eliminate these issues;

however, most remain unresolved. As a result, RFP and freeze casting have a bright

future. However, numerous avenues of research remain open: Parameters must be

developed for building thin wall shells, and interface agents are needed to limit

binder/water interactions. In addition, shell strength and casting surface finishes could be

improved. Finally, pattern loss must be addressed.

To reduce premature cracking during cooling (post pouring) at Mercury Marine,

ceramic strengthening techniques and binder systems (e.g. colloidal silica) were modified

to increase shell toughness. Four-point bend test bars enabled quantitative comparison of

changes made to the shell mold system. Quantitative image analysis allowed for the

comparison of oxidation amounts normalized by blade length.

3

I. Parametric Modeling of the Autoclave De-waxing Process

Edward A. Druschitz

Missouri University of Science and Technology, Rolla, Missouri

Keywords: Investment Casting, Autoclave, Shell Cracking

4

1. ABSTRACT

Snow [1998] suggested that up to 90% of all shell cracking is a result of the

autoclave dewaxing cycle. The majority of cracks are caused by bulk expansion of the

wax as it is heated. The expanding wax stresses the shell, and if the stress intensity

becomes greater than the shell’s strength, the shell cracks. Minimizing the expansion of

the wax during melting eliminates these cracks. A clearer understanding of the factors

affecting the melt profile of the wax can be gained by using computational fluid

dynamics (CFD) to model the interaction among 1) the thermal conductivity of the wax,

2) the thermal conductivity of the shell, and 3) the temperature of the autoclave. The

most favorable melt profile results from a high autoclave temperature (438°K to 458°K)

and saturated thermal conductivity of the shell (1.36 to 1.40 Wm-1k-1) in conjunction with

a low wax thermal conductivity (0.33 Wm-1k-1). These parameters reduce the likelihood

of shell cracking as a result of wax bulk expansion.

2. INTRODUCTION

Investment casting foundries use a saturated steam autoclave to remove wax

patterns from ceramic shells made of fused silica. Cracks nucleated by the autoclave may

result in leakers (run-outs), dimensional distortion, surface defects, and inclusions. The

majority of these cracks are caused by the bulk expansion of the wax during melting.

Minimization of this expansion would drastically reduce shell cracking and associated

defects. Figure 2.1 illustrates the autoclave de-waxing cycle.

5

The autoclave process occurs in the following steps:

1. The shells are placed inside the autoclave.

2. Water is flash boiled into steam and injected into the autoclave.

3. Steam condenses on surfaces of sub-superheated steam temperatures.

4. Water permeates the shells (increasing thermal conductivity).

5. Wax begins to melt at the shell/wax interface causing a volumetric

expansion.

Cracking is a nucleation and growth process. Cracks form in the autoclave when

stresses caused by the bulk expansion of the wax exceed the strength of the shell. Cracks

grow along the path of least resistance in order to alleviate the stress caused by bulk

expansion. Stresses can be relieved by the lateral flow of liquefied wax through vents

and gating.

6

Figure 2.1. Schematic of heat transfer in the autoclave. Steam condenses on the shell’s surface and saturates the shell. Heat is transferred through the shell to the under

lying wax, resulting in melting [Gebelin, 2001].

Snow [1998] estimated that a dry fused silica shell has a thermal conductivity of

2.00 X 10-5 BTUsec-1in°F (1.5 Wm-1K-1). He also estimated that a shell with 25%

porosity filled with water would have a conductivity of 4.21 X 10-5 BTUsec-1in°F (3.15

Wm-1K-1). Kruse and Richards [2005B] measured the dry shell’s thermal conductivity at

0.5 Wm-1K-1 and the saturated shell’s thermal conductivity at 1.4Wm-1K-1.

Once the shell surface is saturated with water, the pressure inside the autoclave

puts the shell in a compressive state of stress until the wax expands volumetrically.

Fused silica undergoes negligible expansion as it is heated from room temperature to the

autoclave operating temperature.

In early instrumented autoclave trials, Jones et al. [2001] found that the interior of

an autoclave reaches its maximum temperature and pressure in less than ten seconds.

7

Later work conducted by Kruse and Richards [2005A] showed that within forty seconds

the outer surface of the shell reached ambient autoclave temperature. In either case,

during the de-waxing process the environmental temperature boundary condition occurs

quickly at the surface of the shell. This is considered in modeling in that, the rapid

development of this condition allows the modeler to ignore the remainder of the

autoclave and apply a constant temperature boundary condition at the shell surface when

constructing the model.

Mathematical models for estimating the thermal conductivity of ceramic-water

porous phase composite structures include: Maxwell, Sson–Frey, Russel, and Bruggeman

[Kruse and Richards 2005B]. These methods result in a dry shell thermal conductivity

range of 0.05 to 0.20 Wm-1K-1 and a water saturated thermal conductivity range of 0.20 to

0.80 Wm-1K-1 [Kruse and Richards, 2005B]. Kruse and Richards [2005B] determined

that dry and water saturated shell thermal conductivities are 0.5 and 1.4 Wm-1K-1,

respectively. They interpreted the discrepancy between modeled conductivity and

measured results stemmed from the inability of previous models to account for a

combination of continuous and non-continuous phases within the shell. Snow [1998]

assumed condensed water from the steam was pulled into the shell via capillary action

due to high pressures. Kruse [2005] and Kruse and Richards [2005B] proposed a

modified Maxwell model that accounted for the change in the thermal conductivity of a

continuous phase as a function of temperature. This model best fit the experimentally

measured data.

These data were used to generate a parametric model to determine the heat flow

through the shell and melt front progression through the wax. The wax pattern’s thermal

8

absorption consists of a sensible heat increase, Cp∆T, and a latent heat of fusion during

melting, mliq∆Hf [Kruse and Richards, 2005B].

The main focus of the work presented here is the development of a parametric

thermal model to evaluate for the influence of 1) pattern wax thermal conductivity, 2)

investment shell thermal conductivity, and 3) autoclave temperature on melt front

progression and its resulting effects on the bulk expansion of the wax.

3. BOUNDARY CONDITIONS AND PARAMETERS

A rectangular block of wax (10 mm×100 mm×100 mm) was chosen for the

current study to allow for an infinite plate approximation and to reduce the edge effects.

The shell was set at 12 mm thick (the average of shells from three industrial sources).

The shell surrounded the entire wax block, as shown in Figure 3.1. Heat transferred

through the shell to the underlying wax resulted in melting and a thermal gradient

through the wax.

9

Figure 3.1. Wax block surrounded by ceramic investment casting shell with boundary

condition Tw (temperature at wall).

The thermal resistance between the materials at their boundaries was ignored

because the thermal conductivity of the shell material is low. Fluid flow was also ignored

for the purpose of this parametric model. A semi-infinite plate solution boundary

condition allowed the edge effects to be ignored. This permitted simplification of a three

dimensional computer model to a two dimensional simulation.

For this parametric model, the fixed boundary condition was a constant shell

surface temperature (Tw) set at values of 433oK, 438oK, and 458oK (Tw) as per the results

of Jones et al. [2004] work on the thermal profiles of autoclaves. Table 3.1 lists the

conditions and variables used for ten simulation runs. The three variables in these tests

were: a) wax thermal conductivity, b) shell thermal conductivity, and c) autoclave (shell

surface) temperature.

10

Table 3.1. Parameters used to determine the influence of shell and wax conductivity and autoclave temperature on the melt profiles of wax during autoclaving.

Name

Shell Thermal Conductivity

(Wm-1K-1)

Wax Thermal Conductivity

(Wm-1K-1)

Autoclave Temperature

(°K)

Model 1 0.55 0.33 438

Model 2 0.55 0.5 438

Model 3 1.36 0.33 438

Model 4 1.36 0.5 438

Model 5 1.4 0.33 438

Model 6 1.4 0.5 438

Model 7 1.36 0.33 433

Model 8 1.36 0.5 433

Model 9 1.36 0.33 458

Model 10 1.36 0.5 458

The shell’s thermal conductivity was varied between three different values: 0.55,

1.36, and 1.4 Wm-1K-1, representing a “dry” (model one and two) and a fully water

saturated shell respectively. The last two values (1.36 and 1.4) were chosen to determine

the sensitivity of the modeled system to small changes in wax thermal conductivity. The

typical autoclave cycle is 30 to 40 minutes. Backup coats saturate 15 seconds after the

door is closed. The primary coat saturates in 80 seconds. Therefore, saturated shell

conductivity was used in the remaining simulations. Wax thermal conductivity was

varied between 0.33 Wm-1K-1 (for low density polyvinyl ether polymer) and 0.5 Wm-1K-1

(for high density polyvinyl ether polymer). Additional properties required for the

simulation are summarized in Table 3.2; the additional properties of the low density

polyvinyl wax were held constant regardless of thermal conductivity to eliminate their

impact on the results.

11

Table 3.2. Additional properties of the wax and shell required for the completion of the parametric simulation.

A quadrilateral face with a three-node edge and both tetrahedral and cubic

volumes was used to mesh the shell and wax respectively. All calculations were

conducted assuming an unsteady state with segregated calculations. Each time step was

recorded in Fluent at 2.0 seconds, with a computational time step of 0.01 seconds. At

each time step, the thickness of the melt front was calculated and saved in an Excel

spreadsheet. Complete melting was defined as a liquid fraction greater than 85%. Figure

3.2 is an example of the 2D slice generated by Fluent software and used to predict the

thickness of the wax melt front. In this figure, blue indicates solid material and red

indicates melted material. Melt front thickness was determined by measuring the

thickness of the red area at a cross-section of the plate’s center. Heat was transferred

through the shell to the underlying wax, causing it to melt. This transfer resulted in a

thermal gradient due to low thermal conductivity of the wax.

12

Figure 3.2. Melt profiles of the wax as a result of the calculated distribution of the liquid fraction at 480s (left) and 640s (right).

4. COMPUTATIONAL PROCEDURES

The enthalpy-porosity technique was used to model the phase change process.

The melt interface was not tracked explicitly; instead, the liquid fraction associated with

each control volume in the domain and computed each iteration. The liquid fraction,

therefore, varied between zero (solid) and one (liquid). The energy equation is written in

terms of sensible enthalpy, h, defined as equation 1

ref

T

ref pTh h c dT= + ∫ (1)

where href in J is the reference enthalpy, Tref in K is the reference temperature, and cp is

specific heat at constant pressure in Jkg-1K-1, and is a function of temperature T. The

enthalpy can be computed as the sum of the sensible enthalpy h and the latent heat ΔH

(equation 2)

Shell

Melted Wax

Un-melted Wax

13

HhH Δ+= (2)

In addition, the latent heat content ( HΔ ) may vary between zero (solid) and L (liquid),

the latent heat of the material. As a result, the liquid fraction (β ) can be defined as

equation 3 if TSolidus≤T≤TLiqudus.

solidusliquidus

solidus

TTT

LH

−−

=Δ

=β (3)

For phase change problems, the energy equation is written as equation 4

( ) ( ) ( ) ( ) STkx

hux

Ht

ht i

ii

+∇∂∂

=∂∂

+Δ∂∂

+∂∂ ρρρ (4)

where H, h , and HΔ in (J) is the enthalpy of the wax, ρ (kgm-3) is the density of the wax,

k (Wm-1K-1) is the thermal conductivity, T (K) is the temperature, and S is the source

term, t is the time in seconds, ui is the fluid velocity in ms-1, and xi is Cartesian coordinate

directions. Fluid flow is irrelevant, thus the velocity term ( )hux i

i

ρ∂∂ is reduced to zero.

Using Equation 3, the sensible enthalpy (h) and the latent heat content ( HΔ ) equal the

enthalpy (H). As such, equation 4 can be reduced to Equation 5.

( ) ( )i

H k Tt xρ∂ ∂

= ∇∂ ∂

(5)

14

This leaves five unknown variables: H, T, h, ΔH, β and five equations (Eq.1-3 and 5).

5. RESULTS

Reporting the results requires the definition of several critical times:

Tms: the time required for melting to begin on the outer surface of the wax.

Tmo: the time required for melting to finish on the outer surface.

Tmc: the time required for melting to begin in the center of the wax.

Tmi: the time required for melting to finish half way to the center.

Tmf: the time required for melting to finish at the center of the sample.

DT: the time difference between the completion of melting on the outer wax surface and

the beginning of melting at the center.

The results of all simulations are shown in Table 5.1. Low conductivity shells

resulted in negative DT times (-72 and -134). Saturated shell conductivity resulted in a

positive DT time. Increased autoclave temperature, saturated shell conductivity, and low

wax thermal conductivity resulted in the highest DT time of 48 seconds.

Table 5.1. Times required for melting given varying shell and wax conductivities and autoclave temperatures. Increased autoclave temperature, saturated shell conductivity,

and low wax thermal conductivity resulted in the highest DT time of 48 seconds.

Tms Tmo Tmc Tmi Tmf DT1 0.55 0.33 438 154 352 280 662 878 -722 0.55 0.5 438 156 374 240 648 802 -1343 1.36 0.33 438 60 138 178 312 484 404 1.36 0.50 438 62 138 144 286 412 65 1.40 0.33 438 58 134 174 306 476 406 1.40 0.50 438 60 134 142 278 404 87 1.36 0.33 433 62 144 180 328 504 368 1.36 0.50 433 64 144 146 300 430 29 1.36 0.33 458 54 118 166 266 416 48

10 1.36 0.50 458 54 118 134 242 354 16

Elapsed Time to Event (sec)Model

Shell Thermal Conducivity (W/(m*K))

Wax Thermal Conducivity (W/(m*K))

Autoclave Temperature (K)

15

6. DISCUSSION

A high DT value will reduce the stress on the shell caused by the bulk expansion

of the remaining wax. Therefore, the outer layer of the wax should ideally finish melting

(Tmo) prior to significant temperature increase at the center of the pattern (Tmc).

The first set of simulations (model one and two) studied the effect of the waxes’

thermal conductivity on melt times. When the thermal conductivity of the shells and the

autoclave temperature were held constant at 0.55 Wm-1K-1 (dry) and 438°K respectively,

the low conductivity of the shell limited the amount of heat transferred to the wax. In

both cases, the wax began to melt at its center before it finished melting at its surface.

The low thermal conductivity of the shell is unfavorable since it will lead to a large bulk

expansion (indicated by a negative DT time).

Models three through six demonstrate that saturated shell conductivities lessen the

effect of wax thermal conductivity on the time required for melting to begin and finish on

the surface (Tms and Tmo). However, the higher thermal conductivity wax (models four

and six) began to melt in the center six seconds after it finished melting at the surface.

The higher shell thermal conductivity in conjunction with lower conductivity wax

(models three and five) result in a more favorable melt profile (larger DT), reducing the

likelihood that the wax will undergo bulk expansion and result in shell cracking.

The last of the simulations (models seven through ten) determined the impact of

autoclave temperature. Wax thermal conductivity did not affect the time required for

initiation and completion of melting on the surface. The increased autoclave temperature

and low conductivity wax resulted in the greatest DT time of 48 seconds.

16

7. CONCLUSIONS

Favorable profiles are defined as having the greatest possible time delay between

completion of melting on the wax’s surface and beginning of melting at the wax’s center

(DT). These profiles should limit the bulk expansion of the wax and thereby reduce the

likelihood of shell cracking. Low thermal conductivity dry shells (0.55 Wm-1K-1) in

conjunction with high wax conductivity (0.5 Wm-1K-1) produced the least favorable

melting profiles (DT of -134 seconds). High autoclave temperatures (438-458°K), high

conductivity saturated shells (1.36-1.40 Wm-1K-1), and low conductivity wax (0.33 Wm-

1K-1) resulted in the most favorable melt profiles (DT of 48 seconds).

8. FUTURE WORK

Future work should include the addition of wax expansion data in order to

calculate the stress applied to the shell. Data on the flow of fluid out of the shell should

also be added to the simulation in order to calculate the alleviation of stress.

17

9. REFERENCES

American Foundrymen’s Society. Handbook on the Investment Casting Process. Des Plaines, Illinois: American Foundrymen’s Society, 1993.

Gebelin, J. & Jones, S. “Modeling of the De-Waxing of Investment Cast Shells”. TMS 2001.

Jones, S. Jolly, M. Blackburn, S. Gebelin, J. Cendrowicz, A. and Lewis, K.

“Measurements of autoclave thermal profiles during high pressure steam de-waxing of investment shells: Part 1 – Vessel profiles.” Materials Science and Technology, May 2005. Vol. 20.

Jones, S. Jolly, M. Gebelin, A. Cendrowicz, A. & Lewis, K. “To Boldly Go Where No

Woman Has Gone Before: Dewaxing Results From FOCAST.” ICI 49th Annual Meeting, 2001.

Kruse, B. “Mold and Metal Interactions in Highly Alloyed Steels”, M.S. thesis,

University of Missouri-Rolla, 2005. Kruse, B. & Richards, V. “Success of a Data Acquisition System Designed to Measure

Thermal, Moisture and Pressure Profiles in Production Autoclaves”. ICI 53rd Annual meeting, Paper #15, 2005.

Kruse, B. & Richards, V. “Thermal and Moisture Characterization During Autoclave

Dewaxing in Investment Casting.” SFSA T&O Conference, Paper # 5.5, 2005. Snow, J. “What Happens During Autoclave Dewaxing”. Investment Casting Institute

46th Annual Technical Meeting. 1998.

18

II. Investment Shell Building on Ice Patterns

Edward A. Druschitz


Keywords: Investment Casting, Rapid Freeze Prototyping, Ice Casting

19

1. ABSTRACT

This work developed shell building techniques for rapid freeze prototyping (RFP). Thin

wall ice patterns were invested by coating patterns with a tridecane interface agent, using

a 20 wt% ethyl silicate binder with 20 wt% fiber-containing fused silica flour added after

the primary coat, delaying the application of the catalyst by 4 hours, and maintaining a -

10°C environment. Finished shells were quantitatively evaluated by determining mold

cavity dimensional reproducibility and shell strength. Tridecane proved an effective

interface agent and resulted in stronger surface coats. It was particularly beneficial when

combined with greater pattern thermal mass, which delays melting for a longer period of

time. A -10°C or lower environment increased shell strength and surface finish.

Tridecane resulted in a 7% loss of thickness in ice 3.175mm (0.125 inches) and thicker.

2. INTRODUCTION

Water does not undergo volumetric expansion as it melts. Yodice [1991, 1998,

1999] first proposed water (ice) as a pattern material for investment casting. Rapid freeze

prototyping is a solid freeform fabrication technique wherein water droplets form frozen

layers of ice that generate thin-wall investment casting patterns. This method and

material combination allows investment casters to produce prototype patterns rapidly,

and it obviates the disadvantages of previous technologies.

Wax undergoes a volumetric expansion during melting, which can result in shell

stresses great enough to cause shell cracking. Commercially available rapid prototyping

polymers undergo a greater volumetric expansion than typical pattern waxes, increasing

the likelihood of shell cracking.

20

Rapid prototyping, or solid freeform fabrication (SFF), is a commercialized

method for quickly producing three-dimensional parts via layer-by-layer deposition. This

technique allows manufacturers to produce prototype parts rapidly, while decreasing

development time and reducing cost, and increasing quality. Typical rapid prototyping

materials undergo greater volumetric expansion than standard pattern wax or demonstrate

such poor fluidity compared to industrial pattern waxes that additional drain offs must be

cut into the ceramic mold, adding man hours and reducing overall part quality.

Figure 2.1. is an image of RFP equipment; the XY table moves, allowing the

nozzle to deposit water droplets on the liquid nitrogen chilled substrate. A freezer houses

the entire unit; water is pumped to the nozzle via an external pump. Continuing

development of this process for industrial use focused on taking thin-walled ice patterns,

building investment shells on them, determining ice pattern thickness reproducibility and

producing metal-matrix-composite (MMC) aluminum castings.

Figure 2.1. Rapid freeze prototyping (RFP) machine inside a deep freezer. Water is deposited on the liquid nitrogen chilled substrate, which is moved by the XY table.

Elevator

Freezer

Nozzle XY Table

Substrate

Pipe

Pump

Motor Driver

21

Can molding involves pouring slurry around a pattern contained in a rigid

structure (i.e., a can). After the slurry hardens, the pattern is melted out and metal is cast

into the remaining void. Jose’s [2005] research applied can molding and casting to

threaded test articles and dental fixtures for dimensional analysis using ice patterns. He

maintained a constant freezer temperature of -16°C. Water based colloidal silica binder

systems could not be used to prevent ice pattern melting, so an ethyl silicate binder was

chosen. Alumino-silicate flours and a triethanolamine catalyst were used to create

investment casting molds. Particle size was 0.075 mm for the alumino-silicate flour,

which was dried at 100°C for one hour before use. Ten weight percent ethyl silicate

binder was diluted by 50% with ethanol to improve moldability. Jose found 46% solids

loading was optimal for can molds. Jose [2005] used Grey Matter, a commercial

alumino-silicate flour with small inorganic glass fibers premixed to increase strength and

improve resistance to cracking during layered shell building.

Investment casters use mechanical mixing slurry tanks rotating at 15 to 18 rpm to

prevent settling of the slurry. Past studies at Missouri University of Science and

Technology noted that mixing slurry before dipping resulted in an increase in slurry

temperatures of up to 10°C due to particle friction. This increase promoted pattern loss

due to melting. A slurry tank was designed and built to fit inside a freezer, allowing the

slurry to maintain proper suspension (i.e., preventing settling) and temperature (-15°C).

The initial design of the slurry tank system is shown in Figure 2.2. A five gallon

bucket was rotated at 16 RPM. A cover reduced alcohol evaporation by increasing the

local vapor pressure a K-type thermocouple monitored slurry temperature. An image of

the system at work inside the freezer is also shown in Figure 2.2.

22

Figure 2.2. Schematic of slurry tank system. It is driven by an electric motor with built-

in reducing gears (16rpm) rotating a five gallon bucket using a ½ inch drive belt.

A step plate pattern allowed for quantitative analysis of shell quality (surface

finish) and pattern loss. The step plate’s length and width allowed each step to be viewed

as a semi-infinite plate. Step thickness varied between 0.38 mm (0.015 inches) and 6.35

mm (0.25 inches) thick, as shown in Figure 2.3.

23

Figure 2.3. Step plate test article used for ice pattern, thin wall, investment casting trials

(units are inches).

The shell was built in seven layers (Figure 2.4). The first layer (primary coat)

determined the casting’s surface finish and quality. The second and third layers (detail

layers) build part detail. The fourth, fifth, and sixth layers (backup layers) produced the

shell’s strength. The final layer (a seal coat) was not stuccoed, and serves to bind the

previous layer of stucco.

Figure 2.4. Diagram of shell coats: The primary coat determines surface finish, detail

coats build pattern detail, backup coats provide strength, and the seal coat holds the last backup layer’s stucco in place.

Ice Pattern

Primary Coat

Detail Coat

Backup Coat

Seal Coat

24

3. EXPERIMENTAL PROCEDURE

The interaction between pattern melt off (water) during the shell building process

dilutes the primary coat, reducing the silica chain length and weakens the primary coat.

Delaying catalyst application allows the ethanol carrier to evaporate, resulting in less

dilution. Industry best practices delay catalyst application by 0.75 to 1.5 hours at room

temperature for wax patterns; allowing ethanol to evaporate.

3.1. PATTERN LOSS

The goal of the pattern loss test was to determine if tridecane would reduce the

interaction between the pattern and the slurry, thus reducing pattern loss. Twelve ice

step-plates were produced for this test. Six of the plates received a double coating of

tridecane; applied by dipping. The other six received no tridecane coating.

The primary coat consisted of 6900 mL of 10 wt.% ethyl silicate and 17

kilograms of a 50/50 mixture of alumino-silicate and fused silica flour. Remaining layers

included 10% fiber to reinforce the fused silica flour. Primary coats did not contain

fibers due to their potentially detrimental effect on surface quality. Step plate patterns

were dipped in slurry, stuccoed with alumino-silicate sand, sprayed with catalyst (a 50/50

mixture of triethanolamine and ethanol), and cured for six hours. Stuccoing sand varied

according to coat. Primary coat stucco was alumino-silicate 0-15 percent in sieve 150,

70-86 percent USS sieve 100, and 5-10 percent in sieve 50. Detail stucco was alumino-

silicate of 9-22 percent sieve 100, 30-44 percent sieve 70, and 30-48 percent sieve 50.

Backup stucco was alumino-silicate 27-37 percent sieve 40, 32-47 percent sieve 30, and

15-25 percent sieve 20. The seal coat was not stuccoed.

25

Images were taken of ice patterns prior to and after shell building and dewatering.

Image analysis software was used to determine the initial thickness of the ice pattern

steps and the final thickness of the shell cavity. Images were converted to grayscale for

thresholding and adjusted so only the ice pattern (before) or shell cavity (after) was

visible. Height measurements taken in 50 pixel increments were used to determine

pattern loss as a function of initial thickness.

3.2. DELAYED CATALYST APPLICATION

A second test determined how long to delay catalyst application to maximize shell

strength. Twelve four-point bend test bars were produced for each of three conditions.

After stuccoing, catalyst was applied to shells with delays of zero, two, and four hours for

each condition respectively and cured for a minimum of six hours. The primary coat

slurry consisted of 17 kilograms of 200 mesh fused silica and 6900 ml of 20 wt% pre-

hydrolyzed ethyl silicate. The remaining layers included fused silica flour, which

contained ten percent fiber by weight. Slurry temperature was maintained at -13°C

during shell building. Stuccoing was performed as described in section 3.1. The test bars

were strengthened at 800°C for two hours. Test dimensions were approximately 3.8

inches (96.5 mm) long by 0.7 inches (18.5 mm) wide. Samples were tested using a

Simpson Universal Sand Testing Machine and a four-point bend testing fixture. Load-at-

failure was recorded in Newtons. Four-point bend strength was determined using

Equation 1 [Baratta, 1982]:

243bdPLS = (1)

26

where P = force, L = distance between supports, d = sample thickness, and b = sample

width. This equation holds true only if the wedge stress is negligible between the support

points and the loading points. This was achieved by minimizing the distance between the

loading and support points in relation to sample thickness (typically a ratio of 1.2 to 1.4)

and the almost negligible deflection of the bar before failure.

4. RESULTS

4.1. PATTERN LOSS

Three ice patterns (eight, nine, and twelve), which had not received a tridecane

coating, broke during shelling. Samples one through six (tridecane coated) and samples

seven, ten, and eleven (uncoated) survived shell building and dewatering. Figure 4.1

compares ice pattern starting thicknesses and resulting cavity thicknesses; Figure 4.2

illustrates the percentage loss for each step. Step one was the thickest step; thickness

decreased for steps two and three. Step three was unquantifiable in samples four and five

due to taper. The average coated step thickness of the ice was 7.56mm for step one,

4.87mm for step two, and 2.83mm for step three. Resulting cavity thickness was 7.00mm

for step one, 4.08mm for step two, and 2.56mm for step three. A net loss of seven,

sixteen, and five percent was found for each step respectively.

The average uncoated step thickness was 6.80mm for step one, 4.27mm for step

two, and 1.83mm for step three. Cavity thickness for uncoated plates was 6.85mm for

step one, 4.13mm for step two, and 2.43mm for step three. This resulted in a three

percent net gain for step one, a three percent net loss for step two, and a 43% gain in

thickness for step three. Careful examination of the mold cavities showed the

27

delamination of the primary coat in the third step of the uncoated plates, accounting for

their increased size over the original pattern as shown in figure 4.3.

0

1

2

3

4

5

6

7

8

9

10

1 2 3 4 5 6 7 10 11

Aver

age Tr

idec

ane

Averag

e with

out

Sample

Heig

ht

of

Ste

p (

mm

)

Ice step 1 Shell step 1

Ice step 2 Shell step 2

Ice step 3 Shell Step 3

Figure 4.1. Ice pattern loss and resulting shell cavity thicknesses. Samples one through

six were tridecane coated, samples seven, ten, and eleven were not coated. Delamination of the primary coat caused cavities to show a net increase in thickness.

28

-120%

-100%

-80%

-60%

-40%

-20%

0%

20%

40%

60%

1 2 3 4 5 6 7 10 11

AVG T

ridec

ane

Aver

age Tr

idec

ane

Sample

Pe

rce

nt

Patt

ern

Lo

ss

Step 1 "0.25 inches" 0.635cmStep 2 "0.125 inches" 0.3175 cmStep 3 "0.06 inches" 0.1524 cm

Figure 4.2. Percentage loss based on step thickness. Samples one through six (tridecane

coated) experienced minimal loss of thickness compared to non-coated step plates.

Figure 4.3. Investment shell exhibiting primary coat delamination.

A sample of a tridecane coated pattern shell and a non-coated pattern shell were

examined. Figure 4.4 shows complete primary coat loss in sections thinner than 1.52 mm

(0.06 inches) for both conditions. The thickest two steps of the tridecane coated patterns

retained the entire primary coat.

Delamination

29

Figure 4.4. Step plate shells with and without tridecane coating. Tridecane coating

increased retention of the primary coat. Step plate thickness increases from left to right.


Initially, twelve four-point bend test bars were produced for each delay time.

Several samples broke during sample preparation. Since the ice patterns were not coated

with tridecane, this was presumably caused by the interaction of ethanol and ice. Any

samples that broke during preparation or outside test fixtures inner supports were

considered invalid. Table 4.1 contains the results for zero, two, and four hour delayed

catalyst application times. Each column lists bar number and calculated strength

(N/cm2). There is an 88% probability that delaying catalyst application for four hours

produced stronger shells (16.4 N/cm2) compared to zero and two hours (11.7 and 14.5

N/cm2 respectively).

Table 4.1. Four-point bend test results for delayed catalyst application. A four hour delay resulted in the highest average strength (16.4 N/cm2).

Time Delay One Two Three Four Five Six Seven Eight Average STDEV0 - Hours 10.1 18.2 9.8 10.9 11.4 9.9 11.7 3.22 - Hours 18.3 10.0 22.3 11.1 10.6 14.5 5.54 - Hours 12.3 10.9 9.8 28.0 10.9 15.7 20.3 23.5 16.4 6.8

Sample Number and Strength (N/cm^2)

Solid primary

Coat

Delamination of

Primary Coat

30

5. DISCUSSION

5.1. PATTERN LOSS

Tridecane was not effective in preventing ice pattern loss for thicknesses of 1.52

mm (0.06 inches) and smaller. Thicker steps lost an average of seven percent of their

starting thickness. Tridecane resulted in improved primary coat retention for the two

thickest steps. Greater thermal mass increased the resistance to melting; presumably,

thicker sections require greater heat input to cause melting and were, therefore, less

affected by time outside the freezer during shell building. Increasing ice thickness

required increased heat input to induce melting. Ice at -15°C and 0.40 mm thick was

calculated to begin melting after 31 seconds at room temperature (25°C), whereas ice at -

15°C and 6.35 mm thick will not begin melting for 360 seconds. Increasing the time

delay before melting occurs would reduce pattern loss. The likelihood of water

interacting with the primary coat would also be reduced, thereby increasing the primary

coat strength.


By delaying catalyst application for four hours after stuccoing, shell strength was

increased from 11.7 N/cm2 to 16.4 N/cm2. Providing time for ethanol to evaporate before

catalyst application resulted in higher silica chain lengths and increased strength.

6. CONCLUSION

Shells built on ice patterns suitable for counter gravity casting of metal-matrix

aluminum composites can be produced using:

31

• tridecane interface agent coating

• 20 wt% ethyl silicate binder

• 20 wt% fiber-containing fused silica flour added after the primary coat

• Delay catalyst application by four hours

• -10°C environment.

Tridecane is an effective interface agent that produced stronger surface coats,

particularly when combined with greater pattern thermal mass, which delays melting for

a longer period of time. A -10°C or lower environment increased shell strength and

improved surface finish. Tridecane resulted in reproducible minimization of ice pattern

loss in ice 3.175mm (0.125 inches) thick.

7. FUTURE WORK

Future work should be conducted inside a freezer because ice pattern melting

reduces pattern accuracy and shell strength. Additional methods of tridecane application

should be explored. Further, emphasis should be placed on using RFP shells to produce

actual castings. Finally, the pattern loss versus starting thickness experiment should be

duplicated on rapid prototyped parts.

32

8. REFERENCES

Baratta, F. “Requirements for Flexure Testing of Brittle Materials.” AMMRC TR 82-20, Army Materials and Mechanics Research Center, Watertown, MA, 1982.

Jose, H. “Investment Casting Using Ice Patterns: Solid Mold and Shell Mold Methods”

University of Missouri-Rolla: Thesis, 2005. Yodice, A. “Freeze cast process ready for licensing”, INCAST: International Magazine of the Investment Casting Institute, 11(12), 19-21, 1998. Yodice, A. “Freeze cast process”, US patent 5,072,770,1991. Yodice, A. “Freeze process cuts casting costs”, Advanced Materials and Processes, 155 (4), 35-36, 1999.

33

III. Oxidation During Solidification of 15.5 PH Marine Propellers

Edward A. Druschitz


Keywords: Investment Casting, Oxidation Formation, Shell Strengthening

34

1. ABSTRACT

Oxidation along the leading edge of cast Fe-15Cr-4.5Ni-3Cu (15-5 PH) stainless

steel marine propellers requires costly non-value added finishing. The addition of an

extra seal coat of slurry after autoclaving and a slower cooling rate can reduce this

oxidation from 1.7 to 1.4mm2mm-1 of oxidation per millimeter of blade length. This

work showed that a seal coat applied after autoclaving re-saturated the shell filling in

surface micro-cracking with slurry, delaying cracking and allowing the casting to cool

adequately before exposure to an oxidizing atmosphere.

2. INTRODUCTION

The goal of this investigation was to reduce the amount and severity of oxidation

along the leading edge of cast Fe-15Cr-4.5Ni-3Cu stainless steel (15-5 PH) marine

propellers (Figure 2.1.) (Unless otherwise noted all chemistries are in weight percent).

Figure 2.1. Oxidation along the leading edge of a cast marine propeller.

According to Sosman [1927], the principal crystalline phases of silica are quartz,

tridymite, crystobalite, and fused silica. When fused silica shells are heated above

1470°C (2678°F) they undergo devitrification to form a high temperature phase,

Severe oxidation along a propellers leading edge

35

crystobalite. Sosman [1927] states that crystobalite undergoes a displacive

transformation when cooled to temperatures of 200-275°C (392-522°F) at atmospheric

pressure. Shells are formulated to develop fine cracking structure during cooling,

resulting in easy removal. Transformation to crystobalite is controlled by mineralizer and

time at temperature; typically sodium aids this transformation.

Lehman et al. [1995] notes that fiber additions date back to adobe, a dried clay

reinforced with straw, to increase strength and toughness. A more modern example is

concrete reinforced with steel rebar. Previous work conducted by Richards and Mascreen

[2002] found fibers increased the strength of 4-point bend samples by causing crack

deflection, wherein crack planes tilt and twist around surrounding grains and fibers.

According to Richardson [2006], brittle cracks propagate in low fracture

toughness materials and result in high flaw sensitivity (low flaw tolerance). More recent

work in fracture toughness, discussed by Richardson [2006], found dispersion of

reinforcement materials of a higher elastic modulus will aid the material in carrying loads

without fracturing.

According to Richardson [2006], cracking in a polycrystalline material can be

broken down into three stages: 1) Stress induced energy is stored within the material, 2)

crack nucleation occurs at the critical load on the largest flaw, and 3) stored energy drives

crack propagation. Failure can be avoided by stress delocalization, accomplished by fiber

reinforcement. Richardson [2006] suggests that the elastic modulus of fibers be two

times that of the matrix. Short or “chopped” fibers of random orientation have gained

wide acceptance within the ceramic industry and are utilized in investment casting

foundries in such products as Grey Matter (fused silica flour containing small inorganic

36

fibers) [Nalco, 2004]. Typically, a fiber length-to-diameter ratio of 8:1 is the minimum

necessary to allow proper modulus transfer from matrix to fiber. Randomly arranged

chopped fibers can cause crack deflection and crack bridging, resulting in increased

fracture toughness [Richardson, 2006].

Crack deflection depends on particle (i.e., grain or fiber) shape. Spherical

particles increase toughness two-fold, whereas a disk can result in three-fold gains in

toughness. Rods result in four-fold toughness improvements [Richardson, 2006].

Volume percentage of reinforcement particles is also important, as shown in Figure 2.2.

which indicates that maximum effectiveness is achieved at 0.5 volume fraction regardless

of shape [Richardson, 2006]. Brittle fiber achieved increased toughness via pull-outs, the

expending of matrix energy resulting in reduced available crack propagation energy

[Richardson, 2006].

Figure 2.2 Effect of particle shape and volume fraction on fracture toughness

[Richardson, 2006].

Ceramic shell building in most investment casting foundries follows similar

procedural steps [AFS, 1993]. Patterns and gating are assembled and dipped into

colloidal silica slurry and stuccoed with coarse refractory particles (typically fused silica).

37

Stucco is applied via “rain fall” sanders or by dipping into a fluidized bed [AFS, 1993].

Shell building typically consists of seven layers (Figure 2.3) [AFS, 1993]. The first layer

(primary coat) determines the casting’s surface finish and quality. The second and third

layers (detail layers) build part detail. The fourth, fifth, and sixth layers (backup layers)

produce the shell’s strength. The final layer (a seal coat) is not stuccoed; it binds the

previous layer of stucco.

Figure 2.3. Diagram of shell coats, the primary coat determines surface finish, detail

coats build pattern detail, backup coats provide strength, and the seal coat holds the last backup layer’s stucco in place.

Shells are dried for 24 hours before autoclaving to remove pattern wax.

Autoclaves use high temperature, high pressure steam at eight atmospheres and 170°C

(338°F) to melt wax patterns. Shells are then fired at 871-1093°C (1600-2000°F) to

increase strength and remove residual organics before pouring in batches of three [AFS,

1993]. Shells are removed from firing three at a time, placed on a refractory brick lined

cooling cart, and poured before placing the next batch. This process is repeated until all

Wax Pattern

Primary Coat

Detail Coat

Backup Coat

Seal Coat

38

shells have been cast. Post pouring, castings are cooled on the same cart on which they

were poured. During cooling, the fused silica shell undergoes a displacive transformation

resulting in cracking, allowing for easy removal of the shell from the thin metal casting.

Oxidation severity depends on composition, temperature, air flow, and exposure

time [Lankford et al., 1985]. Blades on the outside of cart are surrounded by cooler air.

Blades inside the cart cool slower, next to another blade of similar temperature (Figure

2.4). Slower cooling delays crystobalite inversion cracking, reducing exposure of the

casting to air. Faster cooling causes the ceramic shells to crack sooner.

Figure 2.4. Illustration of cooling rate based on position. The inside propeller blades are

surrounded by hot castings and will therefore cool slower.

Thermal images were taken and analyzed of castings poured at 1632°C (2970°F),

which cracked at 428-460°C (802-860°F), as indicated in Figure 2.5 by red spots.

Fast Cool

Fast Cool

Fast

Cool Fast

Cool Slow

Cool

39

Figure 2.5. Thermal image of solidifying casting; areas in red indicate a temperature of

428°C (802°F).

EDS imaging showed that oxidation formed during solidification and cooling of

cast 15-5 PH marine propellers was principally chromium oxide Cr2O3 (light areas) and

silicon oxide SiO2 (dark areas), as shown in Figure 2.6. The images were taken at taken

at 20kV and a working distance of 19mm.

Crack initiation begins here

40

Figure 2.6. EDS image of leading edge oxidation of cast 15-5PH Stainless.


3.1. SHELL STRENGTHENING

Post autoclave seal coatings were tested as a method of delaying cracking onset

during solidification. Six conditions were examined for this study, including baseline

production, which received no seal coating. A seal coat of non-fiber-modified slurry and

four fiber-modified slurry seal coats. De-waxed four blade propeller shells were utilized

for these tests. An individual sample was defined as a single propeller blade. Two

propellers (eight blades) were tested for each of the following conditions:

1. Baseline -- standard production (no modifications)

2. Extra seal coat of slurry (no fibers)

3. Extra seal coat with 0.3 wt% fibers




41

Images of each blade were taken with a high resolution digital camera after shake-

out. Quantitative image analysis of oxidation was conducted using image analysis

software. The leading edge length and oxidation area were determined (see Figures 3.1

and 3.2). These measurements permitted quantitative comparison of oxidation between

blades by normalizing the data resulting in oxidation area per blade length (mm2/mm ±

one standard deviation).

Figure 3.1. Oxidation on leading edge of propeller blade.

Figure 3.2. Image analysis representation of oxidation on leading edge of propeller blade.

42

3.2. COOLING RATE EFFECTS

Eight autoclaved shells were given an extra seal coat and twenty-four hours to dry

before burnout and pouring. These molds were placed in the middle row of the cooling

cart (slow cooling). Eight production (non-modified) shells were also cooled on the

middle of the cooling cart and thirty blades were cooled along the outside of the cooling

cart as a control group. Quantitative image analysis was conducted in an identical

manner to that of section 3.1.

4. RESULTS


Table 4.1 contains the results of the shell strengthening test. Standard production

shells resulted in 3.0 ± 0.6 mm2/mm of oxidation per unit blade length on the leading

edge of the propeller castings. An extra seal coat reduced oxidation per unit length to 2.6

± 0.9mm2mm-1 and 1.1 wt% fiber addition reduced oxidation to 2.6 ± 0.8mm2mm-1.

Table 4.1. Oxidation per unit length of blade (mm2mm-1).

Blade 1 Blade 2 Blade 3 Blade 4 Blade 5 Blade 6 Blade 7 Blade 8 Average STDEV Min MaxProduction 1.99 2.75 3.92 3.89 3.11 2.58 2.96 3.03 3.03 0.64 1.99 3.92Extra Coat 1.63 3.01 3.73 1.36 1.83 3.25 2.98 2.96 2.60 0.86 1.36 3.730.3 wt% 1.97 3.81 5.03 5.26 5.41 2.52 3.55 4.39 3.99 1.27 1.97 5.410.8 wt% 3.51 2.94 1.73 7.27 3.28 1.85 1.83 4.30 3.34 1.84 1.73 7.271.1 wt% 3.24 2.85 3.65 1.98 2.20 1.26 2.08 3.14 2.55 0.80 1.26 3.651.5 wt% 2.30 1.92 2.97 3.01 2.49 3.22 3.71 3.37 2.87 0.59 1.92 3.71

Oxidation per unit of Blade LengthPropeller 1 Propeller 2

Based on Chauvenet’s criterion, a statistical means of assessing data outliers, one

data point was removed from the 0.8 wt% fiber data set, reducing oxidation from 3.3 ±

1.8mm2mm-1 to 2.8 ± 1.0mm2mm-1.

43


The average oxidation for production castings on the outside of the cooling cart is

1.7±0.7mm2mm-1 compared to 1.5 ± 0.5mm2mm-1 for inside cooling. As shown in Table

4.2, propellers that received a seal coat and cooled slower had the least oxidation at 1.4 ±

0.5mm2mm-1 (0.3 mm2mm-1 less oxidation per mm of blade length when compared to

outside cooled production castings). Cooling rate was never measured specifically.

Oxidation per unit blade length was 0.2mm2mm-1 lower for production molds cooled

slowly inside rather than outside cooling.

Table 4.2. Oxidation (mm2mm-1) as a function of position

Average 1.7 1.5 1.4STDV 0.7 0.5 0.5Minimum 0.5 0.8 0.4Maximum 2.9 2.6 2.6Sample Size 32 40 32

Test Conditionmm2/mm of

oxidation Outside InsideInside +Seal

5. DISCUSSION


Because a non-fiber reinforced seal coat reduced oxidation to 2.6mm2mm-1 per

unit blade length compared to 2.6-3.9 mm2mm-1 for fiber additions it did not appear that

fibers aided in strengthening the shells. There is a 72 percent probability that an extra

seal coat reduced the oxidation versus standard production based on recorded data. Also,

the 1.1 percent fiber addition showed a 78 percent probability of improvement compared

to production shells. It appears as though a seal coat applied after autoclaving re-

saturated the shell, filling in surface micro-cracking with slurry and adding thickness.

44

The additional seal coat increased the shells’ overall load-bearing capacity which likely

delayed cracking and allowed the casting to cool adequately before exposure to an

oxidizing atmosphere.


There is a 70 percent probability that castings cooled more slowly inside have

0.2mm2mm-1 less oxidation than quickly cooled outside castings. Presumably, the

decreased oxidation is a result of allowing the casting to cool below 800°C before being

exposed to oxygen. There is also a 95 percent probability that casting of shells receiving

an extra seal coat and cooled slower experienced 0.3mm2mm-1 less oxidation compared

to production castings cooled faster.

6. CONCLUSION

By increasing the shell’s load bearing capacity, toughness, the onset of cracking is

delayed, and castings are able to cool enough to prevent severe oxidation (800°C) before

cracking. These factors reduce the amount of oxidation per unit blade length by as much

as 0.3mm2mm-1.

7. FUTURE WORK

A cooling cart surrounded by refractory or encased by insulating fiber boards,

should slow cooling sufficiently. Inert atmosphere during cooling may also reduce oxide

formation. Modeling could determine optimal rates of cooling. Cooling rate and oxygen

content of the air surrounding the castings should also be measured.

45

8. REFERENCES

American Foundrymen’s Society. Handbook on the Investment Casting Process. Des Plaines, Illinois: American Foundrymen’s Society, 1993.

Lankford, W.T., Jr., Samways, N.L., Craven, R.F., McGannon, H.E., editors, “The

Making, Shaping, and Treating of Steel”, 10th edition, Association of Iron and Steel Engineers, Pittsburgh, PA, 2985.

Lehman, R. L., El-Rahaiby, S. K., Wachtman, J. B. Jr., “Handbook on Continuous Fiber-

Reinforced Ceramic Matrix Composites.” Ceramics Infromation Analysis Center, West Lafayette, IN, 1995.

Nalco Company, “Grey Matter”, Naperville, IL, 2003. Richards, V. L., Mascreen, S., “Thermal Expansion of Investment Casting Pattern Wax”,

AFS Transactions, 2003. Richardson, D. Modern Ceramic Engineering, Taylor and Francis Group, Boca Raton,

FL, 2006. Sosman, R. B., “The Phases of Silica”, New Brunswick, Rutgers University Press, 1965.

46

APPENDIX

Modeling of Heat Transfer through Investment Casting Shells:

Method of Determining Shell Thermal Conductivity

Edward A. Druschitz

Simon Lekakh

Dr. Von Richards

1. ABSTRACT

This work demonstrates a simple and inexpensive method for measuring thermal

conductivity of investment casting shells. Reducing shell cracking during the autoclave-

dewaxing cycle is a goal of all investment casting foundries. However, before cracking

can be reduced, the factors that contribute to crack initiation and propagation must be

established. Shell cracking can lead to many surface defects, including heavy oxidation

resulting in pitting. These discontinuities can be removed only by non value-added

manual labor, increasing the overall cost of a casting. Although extensive work on

modeling the autoclave dewaxing process has been conducted at a number of universities

and by numerous research facilities, most of this work has either relied on assumptions

regarding thermal conductivity of the shells or measured shell thermal conductivity using

expensive and bulky equipment plus, including a production autoclave. Such

measurements require production downtime.

2. INTRODUCTION

The investment casting process is similar to that used in the ancient world:

Modern investment casting foundries use robots to dip wax pattern trees, ceramic

47

materials (such as fused silica) are applied in layers to create shells, and steam autoclaves

(boilerclaves) are used to remove the wax from the shells. The basic investment casting

procedure is shown in Figure 1.

Figure A.1. Basic investment casting procedure.

The single greatest factor determining the quality of an investment casting is the

shell into which the metal is poured. No step in the investment casting process affects

shell quality more than autoclaving. Snow 2 suggests that 90% of all shell cracks

originate in the autoclave; fins, dimensional problems, shell debris, and leakers are all

results of these cracks. The autoclave is a pressure vessel that uses high temperature

steam to aid in the extraction of the wax from the shell. However, this relatively simple

machine is a modern day “black box” for the simple reason that quantifying what goes on

inside an autoclave is difficult, requiring expensive equipment that can withstand

pressures up to 10 bar and temperatures as high as 180°C (Gebelin3). During the

autoclave dewaxing cycle, high temperature steam penetrates the porous ceramic shell

(changing the thermal conductivity of the shell), transferring heat to the wax, and thus

48

causing the wax to melt and run out of the shell cavity (Gebelin J-C4). Work conducted

by Connolly5 suggests that the specific heat of a ceramic shell could be calculated using

the rule of mixtures. This equation, which treats each shell component (stucco and

slurry) as an individual element, is labeled below as Equation 1.

Equation 1. (Connolly5)

...332211 CpFCpFCpFCpShell ++=

where:

CpShell = Specific heat capacity of the investment shell as a whole

F1 = Fractional mass of material 1

Cp1 = Specific heat capacity of material 1

Connolly5 used a differential scanning calorimeter (DSC) and Equation 2 to

determine the specific heat capacity of each shell component, i.e., the primary coat slurry,

primary coat stucco, and so on

Equation 2. (Connolly5)

( )βδδ

δδ

RsR

RS

sRSSR CpCpt

TCpt

TCp −=−=Θ−Θ=ΔΘ

where:

SRΔΘ = Differential heat flow rate

49

SΘ = Heat flow from sample

RΘ = Heat flow from reference

CpS = Sample specific heat capacity

STδ = Temperature change in sample

CpR = Reference specific heat capacity

RTδ = Temperature change in reference

tδ = Change in time

β = Average heating rate

Connaly5 applied the specific heat capacity values generated for the individual

components using the DSC to Equation 1 to calculate the overall specific heat capacity

for the shell. He then compared those values to a measured Cp of the whole shell. The

results were very similar, thus proving that the Rule of Mixtures can be used to calculate

the Cp value of an investment shell. However, this procedure did not consider the effects

of water infiltration or condensation in the shell.

Although Connolly’s5 work explained the speed with which the investment shell

heated up in an autoclave, it failed to address the heat transfer through the shell to the

underlying wax. Jones6 corrected this deficiency by building a shell around a copper

plate of known size and heating it in both wet and dry conditions. For the wet tests, the

shelled plate was submerged in cold and hot water, and the temperature change of the

copper plate over time was recorded. Following this procedure, the shell was removed

and the average shell thickness was determined. Rearranging Equation 3 to find the

50

shell’s thermal conductivity, Jones6 was able to determine the thermal conductivity of the

shell based on the heat flow through it.

Equation 3. (Jones 6)

( )⎟⎟⎟⎟

⎠

⎞

⎜⎜⎜⎜

⎝

⎛

−

⎟⎠⎞

⎜⎝⎛

⎟⎟⎠

⎞⎜⎜⎝

⎛=

flCu

Cu

S

SCupCus TT

tT

AlCM

K δδ

,

Analysis of the data showed that the thermal conductivity of the shell was

approximately 0.4 – 0.55 W/moK in the dry condition and 0.9 – 1.1 W/moK in the wet

condition. This result demonstrated that the thermal conductivity doubled when water

filled the porous ceramic shell (Jones 6). This dramatic increase in thermal conductivity

is favorable since it would create a smaller thermal gradient across the shell, thus

decreasing the depth of the wax melt front and reducing the amount of bulk wax

expansion that might cause the shell to crack.

Sabau7 expanded on the work of Heames and Geiger (1978), Huang et al. (1989)

and Hendricks and Engelhardt (1993) by further developing the theory of packed beds as

a useful model to explain the thermal conductivity of sintered (fired) fused silica shells.

Sabau’s7 research examined two methods to measure the thermal conductivity of the

shells, the hot-wire method (ASTM C1113) and laser-flash (ASTM C714). After

preliminary experimentation, Sabau found that the Laser-Flash method was too sensitive

to the shell’s thickness and yielded inconsistent results; therefore, he settled on a form of

the hot-wire test to determine shell thermal conductivity. This work demonstrated that

the hot wire method can determine shell thermal conductivity and emissivity, which

51

could aid in the modeling of shell thermal properties during cooling. A semi-empirical

correlation for thermal conductivity indicated that the radiative component of the thermal

conductivity can, be expressed as Equation 4 (Sabau7).

Equation 4. (Sabau7)

34)( TdETK pr Β=

where:

Kr = Radiative component of thermal conductivity

E = Factor of order correlated to radiation properties

Β = Stefan-Boltzman constant

dp = Particle size

T = Absolute temperature (K)

Sabau used this equation as a basis to determine emissivity, which he would later

use to determine the heat transfer coefficient (HTC) from the shell to the ambient air.

Until research progressed to this point, thermal conductivity of the shells was largely

based on lab experiments that had yet to be verified by actual instrumented autoclave

trials. Using a specially built high temperature test cell, Kruse8 was able to invest a

copper plate and actually record the temperature increase during a typical autoclave

cycle. This data was used to calculate, for the first time, the thermal conductivity of the

shell during the autoclave process using:

52

Equation 5. (Kruse 8)

( )CuSp TT

XAk

tTMC

−⎟⎠⎞

⎜⎝⎛=

Δ

Using this equation, Kruse 8 calculated the thermal conductivity and pressure of a

foundry's shells to be 0.5 W/moK dry and 1.4 W/moK wet. For dry shells, this trial

yielded the same results as published by Jones6 but it showed that high temperature

pressurized steam resulted in a higher thermal conductivity. According to Kruse8, during

the autoclave cycle, the air is compressed 87.5 percent, and the condensed steam is pulled

into the underlying shell by capillary action. During Kruse’s8 tests, the thermal

conductivity increased by 2.0 – 2.5 times depending on shell structure and composition.

Presumably, this increase occurs because the conductivity of the shell changes with both

saturation and temperature, indicating that existing models of two-phase structures did

not adequately represent the shell’s thermal conductivity. Kruse8 proposed a revised

Maxwell model in which the thermal conductivity of the continuous phase (ceramic and

polymer) was also affected by the moisture content. The underlying physics used to

justify this approach was that the polymer and colloidal silica at the contact points of the

refractory grains absorb water and thus change both the cross-sectional area and the

properties of the contact points. This assumption would be important later when Sabau7

adopted the hot wire method to create an inexpensive and non-intrusive test to determine

the thermal conductivity of shells.

Various mechanisms of heat transfer may be involved during shell bulging in the

investment casting process. The shell, which is a highly porous structure, could transfer

53

heat by: 1) thermal conductivity through a skeleton of solid fused silica particles, 2) by

air conductivity in closed pores, with 3) additional air convection in open interconnected

pores and, finally, 4) radiation during firing and pouring at high temperature. In addition,

wax removal in the autoclave is assisted by water vapor, which could also dramatically

change the rate of heat transfer as a result of the high thermal conductivity of water and

heat of condensation liberation. It is difficult to develop a theory that takes into account

all these possibilities. The experimental methods typically used to measure heat transfer

have been based on steady state measurement techniques. These measurements require

that samples be placed between a heat source and a heat sink, and the temperature

gradient reflects the value of the coefficient of thermal conductivity. Unfortunately,

steady state methods do not accurately reflect real non-steady-state industrial processes,

such as shell dewaxing in an autoclave. Non-steady-state measurement techniques are

more attractive because they can provide data representing combined variables, including

temperature, thermal conductivity, heat capacity, and thermal diffusivity.

Thermal diffusivity is a material’s ability to adjust its temperature to the

surroundings quickly; it is expressed in Equation 6. Heat diffusivity is the ability of the

shell to absorb heat (heat diffusivity = kρCp) (Poirier10). Equation 7 shows the heat flux

into a mold. The rate at which latent heat is evolved is shown in equation 8 (Poirier10).

During the solidification of a casting, the amount of solidified material depends on the

characteristics of the metal’s (Tm, To, p and Hf) and the heat diffusivity of the mold

materials (k, p and Cp) (Poirier10). During the autoclave dewaxing process, liquefaction

occurs. The same information needed to determine solidification times is also required to

54

model the autoclave process; however, the parameters for melting wax are used in place

of those for solidifying metal.

Equation 6: (Poirier 10)

pCk

ρα =


( )00

* TTtCk

q mp

X

−=∫= π

ρ


⎟⎠⎞

⎜⎝⎛=∫

= tMHq f

X δδρ

0


tCkH

TTM Pf

M ρρπ ⎟

⎟⎠

⎞⎜⎜⎝

⎛ −= 0(*2

Two standard methods are used for non-steady-state heat transfer measurements.

The first, called the hot wire method, creates a known value of heat energy inside the

media using a micro heater. A K-type thermocouple is used to determine the temperature

55

curve, and the thermal conductivity (k) can be determined from the temperature versus

time curve. The thermal conductivity of the sample can be derived as follows:

Equation 10: (Carlsaw 11)

( )Atk

QTs

+⎟⎟⎠

⎞⎜⎜⎝

⎛=Δ )ln(*

4π

Equation 11: (Yamasue 12)

...2

4ln 2 ++⎟⎟⎠

⎞⎜⎜⎝

⎛=

w

s

e

S

kk

CrA α

where r is the radius of the wire, Ce is Euler’s constant, and α is the thermal duffusivity

of the sample. Typically, the thermal conductivity of the sample is the slope of the linear

relationship between TΔ and ln(t), calculated as:

Equation 12: (Yamasue 12)

1

)ln(4

−

⎟⎠

⎞⎜⎝

⎛ Δ⎟⎠⎞

⎜⎝⎛=

tTQks δ

δπ

The second method is a transient technique, the laser flash method. This method

measures heat diffusivity and requires additional measurements or assumptions in terms

of the value of specific heat capacity.

56


The present work has developed a novel method for dynamic measurement of

thermal processes in unsteady thin shells and bulk sand media. This method is based on

the generation of a stable and known value for an energy impulse created by passing

direct current through a wire microheater, combined with temperature measurements

inside the media near the heat source. The micro impulse of heat has minimal influence

on existing thermal processes and properties of the sand media. The relaxation time after

the current is turned off is short and therefore permits the reproduction of cyclic

measurements of the thermal properties in rapidly changing conditions. Also, the device

simultaneously measures the absolute temperature of the media.

The microheater was made from 0.38 mm Alomega wire and was approximately

15 mm long. The microheater was welded to thicker wire (0.8 mm) of the same material

to concentrate the heat impulse on the measured space. A type K thermocouple (0.38

mm wire for fast response) was used for temperature measurement. Alumina tube (05

mm in diameter) with four holes was used for the shell of the device. A high resolution,

24-bit data acquisition system and programmable power supply were connected to a PC.

Programming was done with LabView 8 software, which supplied precise

voltage/current/time parameters. A schematic of the device is shown in Figure 2.

57

ProgrammablePower Supply

Sand Media

Data AcquisitionSystem

PC

Micro Heater

Thermocouple

Figure A.2. Method of unsteady thermal conductivity measurements.

The new method was first tested using bulk dry sand. Temperature impulses at

various levels of impulse energy are shown in Figure 3. The measured amplitude of the

temperature impulse was increased with increasing electrical current and heating time.

The amplitude of temperature impulse had minimal variations in sequential

measurements when the same electrical impulse was applied. Full temperature relaxation

time increased from 1 minute for a 1A/3 sec heat impulse to 5 minutes for a 3A/60 sec

heat impulse. Relaxation time refers to the minimum possible time between sequential

measurements. The necessary test cycle can be designed with a programmable power

supply.

58

24.0

24.5

25.0

25.5

26.0

0 1 2 3 4 5 6

Time, min

Tem

pera

ture

, 0 C

1 A 5 sec

20.0

30.0

40.0

50.0

0 2 4 6 8 10 12

Time, min

Tem

pera

ture

, 0 C

3 A 60 sec

a) b)

Figure A.3. Measured temperature in bulk dry sand for a) 1A and 5 sec impulse and b) 3A and 60 sec impulse.

The sensitivity of this new method was tested using sand with varying moisture

contents. The last composition was chosen to evaluate the potential influence of media

electro-conductivity on measurement results. In addition, mixtures of dry sand and

mineral oil were tested to determine the potential influence of water on short circuiting

and heat generation. The applied electrical impulses were 2A for 120 seconds. The

initial measurement data is provided in Figure 4a, and the influence of the moisture and

oil additives on the amplitude of the temperature change is shown in Figure 4b. In both

cases, this method indicated that the temperature increase in the sand media with more

thermally conductive liquid (water or oil) was far less than without. The measurement

technique was not affected by the electrical conductivity of the liquid.

59

20

30

40

50

0 10 20 30 40 50 60 70

Time, min

Tem

pera

ture

, 0 C

Dry sand

+ 6% water

+ 1.5% water

0

4

8

12

16

20

0 2 4 6 8 10

Additions, %

Tem

pera

ture

incr

ease

, 0 C

Sand + waterSand + mineral oil

a) b)

Figure A.4. Influence of moisture and mineral oil additions to dry sand on the amplitude of the temperature increase.

4. COMPUTER MODELING

Fluent software was used to model unsteady heat transfer in 3D media with an

internal energy source (wire). Fluent is a finite element modeling package; for this work,

a 3-node edge, quadrilateral face, and hexahedral volume were used. The semicircle

picture in Figure 5a represents the internal energy source, which was operated in the open

position (2A) and closed position (0A) using a constant resistivity. The thermocouple

was located between the two legs of the semi-circle. The boundary and initial conditions

used in this preliminary modeling ignored the thermal resistance between the materials at

their boundaries because the thermal conductivity of the shell material was already very

low. The principle equation used by Fluent is shown in equation 13 (Fluent 13).

Temperature dependent values of heat capacity Cp are shown in equation 14; they were

applied to the wire heater and shell media without additional thermal resistance (Fluent

13).

60

Equation 13: (Fluent 13)

qxTk

xh

t ii

+⎟⎟⎠

⎞⎜⎜⎝

⎛=

δδ

δδρ

δδ

where: ρ is density, h is enthalpy (which equals dTCT

T pref∫ ), k is thermal conductivity, T

is temperature, and t is time.

Equation 14: (Fluent 13)

MediaMedia

WireWire x

TkxTk ⎟

⎠⎞

⎜⎝⎛=⎟

⎠⎞

⎜⎝⎛

δδ

δδ

The model used published thermal properties of bulk dry sand, and the results

were compared to experimentally measured temperature increases at 2A, as shown in

Figure 5.

0

4

8

12

16

0 100 200 300 400 500 600

Time, sec

Tem

pera

ture

incr

ease

, 0 C

MeasuredCalculated

a) b)

Figure A.5. Computed temperature field (a) and comparison of experimentally measured temperature increase with calculated data for dry bulk sand (b).

61

The calculated temperature increase was similar in media with different known

values of thermal conductivity as shown in Figure 6.

y = 11.291x-1.3571

R2 = 0.9966

0

0.5

1

1.5

2

2.5

3

0 4 8 12 16 20 24 28 32 36

Temperature increase, 0C

K, w

/mK

Figure A.6. Calculated temperature increase and K of materials

5. EXPERIMENTAL RESULTS

Three shell probes were created, and investment shells were built on them.

Sample 1 is a production built shell representing the entire shell building process

(primary coat and secondary coats) at a participating foundry. Samples 2 and 3 are from

different foundries; sample 2 received only primary coats, and sample 3 received only

backing coats. Pictures of these samples are shown in Figure 7.

62

Figure A.7. Completed thermal conductivity probes.

Three shells were tested using different conditions (specified in Table 1). The

shells were initially submerged in dry sand and tested. Tests were then conducted in

steam by isolating the samples in a Kaowool box and using a continuously working

steamer at 60-90°C and normal atmospheric pressure. The samples were also tested

while submerged in water for 20 minutes. The temperature increase (°C) was measured

and the thermal conductivity (k) calculated, results are shown in Table 1.

63

Table A.1. Results of initial round of shell thermal conductivity testing.

Test

conditions

Sample 1 - primary &

backing

Sample 2 – primary

only

Sample 3 – backing

only

T, °C

increase

K,

W/moK

T, °C

increase

K,

W/moK

T, °C

increase

K,

W/moK

Initial shell 17.85 0.23 9.29 0.55 10.33 0.48

In steam 9.8 0.52 6 1.0 5.96 1.01

In water 5.9 1.02 4.77 1.36 5.8 1.05

6. RESULTS AND DISCUSSION

Past work conducted by Brandon Kruse8 at the Missouri University of Science

and Technology to develop a similar shell conductivity test required a large insulating

test chamber to shield the data acquisition system from the intense heat and pressure

created inside the autoclave. During his trials, Kruse coated a copper block with a

production ceramic shell similar to samples 2 and 3 described above. Sample 2

demonstrates the thermal conductivity of the primary coat, and sample 3 demonstrates

that of the backing coat. The present work evaluated each of the shell layers separately to

determine if grain size had an effect on the shell's thermal conductivity. Two

observations are note worthy: first, when dry, samples 2 and 3 showed similar thermal

conductivity (0.55 and 0.48 W/m°K). Second, the thermal conductivity of the primary

sand (sample 2) was 1.36 W/m°K when submerged in water. Prior work conducted by

Kruse8 showed similar results in a production shell (with both a primary coat and backing

coats). During his trial runs, the dry shell had a thermal conductivity of approximately

64

0.64 W/m°K, and at saturation in the autoclave the thermal conductivity was

approximately 1.56 W/m°K using the Maxwell model (Kruse 8). The thermal

conductivity of the finer primary sand was 0.31 W/m°K higher than the backing sand

when saturated, likely because the smaller grain size led to high interconnectivity. Thus

heat was transferred through the backing faster by conduction. Also, the density of the

finer prime coat stucco was higher than that of the coarser backing stucco, and pore size

was greater in the backing stucco. Further, the difference between the saturated results of

the hot wire test and Kruse’s8 tests may be the pressure at which the autoclave operates.

At eight atmospheres of pressure, air in the autoclave would be compressed 87.5% and

the steam would be pulled through the shell by a capillary effect (Kruse 8).

7. CONCLUSIONS AND RECOMENDATIONS

The hotwire method permits accurate measurement of the thermal conductivity of

an investment shell without requiring a production autoclave in which to run tests. Also,

the small probes can easily be shipped to the foundry where the samples can be invested

in the materials of interest to the lab. Thus, there is no need for a lab technician to travel

to each location. Knowing the thermal conductivity of an investment casting shell will

permit more precise modeling of what happens inside the autoclave and thus aid in the

understanding and reduction of shell cracking.

Future work for this project will include the creation of numerous sample probes.

These probes will be sent to participating foundries to be shelled, and then mailed back

for testing. These tests will aid in both the development of a more accurate autoclave

model and clarify how various parameters, including shell slurry, stucco, and grain size,

65

affect the thermal conductivity of shells. In addition, this information will aid in the

further development of the autoclave model.

8. ACKNOWLEDGEMENTS

The authors would like to thank Mercury Marine Corporation and O’Fallon

Casting for their support of this project, and the Mercury Marine Propeller Division for

their time, patience, and dedication to this work.

This report is based upon work supported by the U.S. Department of Energy

under Award No. De-FC36-04GO14230. Any findings, opinions, conclusions, or

recommendations expressed here are those of the authors and do not necessarily reflect

the views of the Department of Energy.

66

9. BIBLIOGRAPHY

1. American Foundrymen’s Society, “Investment Casting Process”, American Foundrymen’s Society, 1993.

2. J. Snow, “What Happens During Autoclave Dewaxing,” Proc. 46th Annual Technical

Meeting of the Investment Casting Institute, 1998, paper 5. 3. Gebelin J-C, Jones S and Jolly M, “Modeling of the De-Waxing of Investment Cast

Shells”, computational Modelling of Materials, Minerals, and Metals Processing, TMS, 2001.

4. Gebelin J-C, Jolly M.R and Jones S. “Process Modelling Research for Investment

Casting,” Proc. 48th Annual Technical Metting of the Investment Casting Institute, 2000.

5. Connolly S, Jones S, Marquis P.M, Ford D.A. “Specific Heat of Investment Casting

Shells” 10th World Conference on Investment Casting, Paper 8. 6. Jones S. Jolly M.R. Blackburn S. Gebelin J-C. Cendrowicz A. and Lewis K. “Effect of

moisture upon mechanical properties of ceramic moulds during high pressure steam dewaxing,” Materials Science and Technology, July 2003, Vol. 19.

7. Sabau A.S. and Viswanathan S, “Thermophysical Properties of Zircon and Fused

Silica-Based Shells for Investment Casting,” AFS Transactions, 2004. 8. Kruse B. and Richards V. “Thermal and Moisture Characterization During Autoclave

Dewaxing in Investment Casting,” SFSA T&O Conference Paper No. 5.5: 2005.

9. Holman J.P. “Experimental Methods for Engineers,” McGraw-Hill. Hightstown, NJ. 1994.

10. Poirier D.R. and Geiger G.H. “Transport Phenomena in Materials Processing”, TMS 1994.

67

11. Carlsaw H. S. and Jaeger J. C. “Conduction of Heat in Solids”, Oxford at the Clarendon Press, 1947.

12. Yamasue E, Masharhiro S, Hiroyuki F and Kazuthiro N. “Nonstationary Hot Wire

Method with Silica-Coated Probe for Measuring Thermal Conductivities of Molten Metals.” Metallurgical and Materials Transactions A, Volume 30A, August 1999.

13. Fluent Incorporated. “Fluent 5 User’s Guide.” Fluent Incorporated, Volume 2, 1998.

68

VITA

Edward Alan Druschitz was born on June 8, 1983 in Michigan. He grew up in

Rochester Hills, Michigan, before moving to Troy, Michigan. In the eighth grade, he

moved to Lynchburg, Virginia where he lived until graduating from Jefferson Forest

High School in 2001. He earned his Bachelors degree in Mechanical Engineering

Technology from Central Washington University in 2005 before following his dream to

study metallurgical engineering technology at Missouri University of Science and

Technology (formerly University of Missouri-Rolla) where he earned his master’s degree

in 2009.

69

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